Article pubs.acs.org/IECR
New Pressure Swing Adsorption Cycle Schedules for Producing HighPurity Oxygen Using Carbon Molecular Sieve Nima Mohammadi, Mohammad I. Hossain, Armin D. Ebner, and James A. Ritter* Department of Chemical Engineering, Swearingen Engineering Center, University of South Carolina, Columbia, South Carolina 29208, United States S Supporting Information *
ABSTRACT: Two 3-bed 5-step and two 4-bed 7-step continuous feed pressure swing adsorption (PSA) cycle schedules were systematically devised and studied via simulation for producing high-purity O2 (≥99.5 vol %) at high recovery using a carbon molecular sieve (CMS). Two binary (O2:Ar) and two ternary (O2:Ar:N2) feed mixtures were investigated to mimic the product produced from a typical O2 PSA system. In addition to feed composition, the effects of light end (LE) versus heavy end (HE) equalization (Eq), two sequential Eq steps versus Eq followed by forced cocurrent depressurization (CoD)/light end pressurization (LEP) steps, feed flow rate (i.e., feed throughput θ), cycle time (tc), Eq step valve coefficient (Cv), and countercurrent depressurization (CnD) step Cv on the process performance were studied in terms of O2 purity (Pur, vol %), O2 recovery (Rec, %), and θ (LSTP, h/kg). In descending order, the best performance from each cycle was Pur = 99.5, Rec = 93.3, and θ = 455.7 for a 4-bed 7-step PSA cycle with LE Eq and forced CoD/LEP (III); Pur = 99.6, Rec = 85.7, and θ = 471.4 for a 4bed 7-step cycle with two LE Eqs (II); Pur = 99.5, Rec = 77.4, and θ = 1089.5 for a 3-bed 5-step cycle with LE Eq (Ia); and Pur = 99.5, Rec = 68.7, and θ = 628.6 for a 3-bed 5-step cycle with HE Eq (Ib). The two Cv values exhibited only marginally unfavorable effects on Pur at smaller values. All four cycles were very sensitive to both tc and θ, with Pur decreasing and Rec increasing with decreasing θ, and Rec decreasing and Pur exhibiting a maximum with increasing tc. For cycle III, the Pur exhibited a maximum with a change in the CoD end pressure, with more negative effects on Pur at longer tc. Changing the levels of Ar or N2 in the feed at fixed O2 content had essentially no effect on Pur or Rec, while decreasing the level of O2 in the feed had an almost proportionally negative effect on Pur and no effect on Rec. When N2 was replaced with Ar in the feed, an unexpected equilibrium effect counteracted an expected kinetic effect, producing a counterintuitive negative result on Pur.
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INTRODUCTION Adsorptive air separation into O2 and N2 enriched products are both major commercial applications for pressure swing adsorption (PSA).1 Most commonly, an enriched O2 product can be produced from air based on an equilibrium-controlled separation using a molecular sieve zeolite such as 5A or LiLSX, or a N2 enriched product can be produced from air based on a kinetically controlled separation using a carbon molecular sieve (CMS). However, the corresponding enriched products both contain Ar because it adsorbs and thus enriches like O2 in molecular sieve zeolites, and it diffuses and thus enriches like N2 in CMSs.2 Ideally, this means that a N2-free O2 product contains about 5 vol % Ar and an O2-free N2 product contains about 1.2 vol % Ar. Because Ar diffuses like N2 in a CMS and adsorbs like O2 in a molecular sieve zeolite, there has been considerable effort to produce a high-purity O2 product (>99 vol %) using various combinations of equilibrium and kinetic separations in singleand two-stage PSA systems; adsorbents that selectively adsorb both Ar and N2 over O2 have also been developed and used in © XXXX American Chemical Society
PSA processes to produce high-purity O2. These two topics have been recently reviewed in detail.3−5 The works of interest to this study include those of Rege and Yang,6 Kim et al.,7 Jee et al.,8 and Jin et al.9 because they developed various CMS-based PSA processes to enrich either O26−8 or Ar6,9 from a feed stream consisting of 95 vol % O2 and 5 vol % Ar or something close to this composition, possibly containing some N2. This feed stream is essentially the O2 product produced from a typical equilibrium-based air separation PSA unit. These four studies6−9 all proposed various PSA cycle schedules, sometimes in terms of the number of beds, cycle steps, and their durations and other times in terms of just the cycle steps and their durations. Some were also clear that the feed step is discontinuous. Others did not address whether the feed step is continuous or discontinuous. Received: July 5, 2016 Revised: September 8, 2016 Accepted: September 15, 2016
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DOI: 10.1021/acs.iecr.6b02570 Ind. Eng. Chem. Res. XXXX, XXX, XXX−XXX
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Industrial & Engineering Chemistry Research Continuous feed is perhaps the most important aspect of any PSA cycle schedule. Although it is certainly possible to operate a PSA process with a discontinuous feed step, it is hardly ever practiced because compressors are not as efficient during intermittent operation. Even if a source of feed is available at the desired feed pressure, which is the case if the feed is being supplied from the product of an equilibrium-based O2 PSA unit, it is still prudent to use all of the O2 product to maximize the recovery of the high-purity O2 product from this train of PSA systems. However, in these four studies6−9 this aspect was either completely overlooked or the cycle schedule was simply stated to have a discontinuous feed step, as mentioned. In the Supporting Information, the cycles utilized in these four studies are described along with their performance, and continuous feed PSA cycle schedules are devised for each one using the minimum number of beds in each case, and for a few cases the maximum number of beds are used to avoid the use of an excessive number of idle (I) steps. It suffices to state that the analysis of the PSA cycle schedules in the literature6−9 presented in Supporting Information produced in some cases unusual continuous feed PSA cycle schedules that never achieved the rather stringent performance required in this work. Moreover, the addition of idle steps to make some of them continuous feed PSA cycle schedules may have detrimental effects on the process performance, especially because they are all based on kinetic separation. Therefore, the objective of this work is to develop a very efficient (high feed throughput) continuous feed PSA cycle schedule for producing a high-purity O2 product (>99.5 vol %) at high recovery (>90%) using a CMS. Based on considerable intuition10−13 and many trials, two 3bed 5-step and two 4-bed 7-step continuous feed PSA cycle schedules were systematically devised and studied via simulation using a dynamic adsorption process simulator (DAPS). This version of DAPS is similar to, but more advanced than that reported on previously.14 The feed to these PSA processes could come from the O2 product of any typical equilibrium-based air separation PSA unit. In this case, it would be obtained from Chart Industry’s Eclipse medical oxygen concentrator that NASA is considering as a source of O2 for use during medical emergencies in space. The CMS PSA unit would continuously use the O2 product from the Eclipse PSA unit to produce 200 lbs/yr of 99.5+ vol % O2 (i.e., containing cycle II > cycle Ia > cycle Ib. The 4-bed 7-step PSA cycle with LE Eq and forced CoD/ LEP (cycle III) resulted in an O2 purity of 99.5 vol % at an O2 recovery of 93.3% and θ = 455.7 LSTP/h/kg. This was followed by the 4-bed 7-step cycle with two LE Eqs (cycle II), which provided an O2 purity of 99.6 vol % at an O2 recovery of 85.7% and θ = 471.4 LSTP/h/kg. These two performances were similar except for recovery. In contrast, the 3-bed 5-step cycle with LE Eq (cycle Ia) provided an O2 purity of 99.5 vol % at θ = 1089.5 LSTP/h/kg but with an O2 recovery of only 77.4%. Although the worst performing PSA cycle was the 3-bed 5-step cycle with HE Eq (cycle Ib), its performance was still reasonably good. It resulted in an O2 purity of 99.5 vol % at an O2 recovery of 68.7% and a θ of 628.6 LSTP/h/kg. Not surprisingly, the two 3bed 5-step PSA cycle schedules exhibited the highest feed throughputs, but the corresponding recoveries were low. Clearly, the additional Eq or CoD/LEP cycle steps that required an additional bed to be added to the 3-bed PSA cycle schedule significantly improved the O2 recovery at the expense of decreasing the feed throughput. In general, the results in Tables S5−S7 show that it was relatively easy to obtain an O2 purity >99.5 vol % at O2 recoveries >50% and at a reasonable θ. However, achieving an O2 purity >99.5 vol % at O2 recoveries >90% and at a reasonable θ was quite challenging. It was achieved in just two
Figure 6. Performance of the 3-bed 5-step PSA cycle schedules Ia and Ib: effect of feed throughput and cycle time on the O2 purity and O2 recovery with (a) light end and (b) heavy end equalization steps. The cycle step sequence is F-EqD-CnD-EqU-LPP.
3-bed 5-step PSA cycle schedules (cycles Ia and Ib). The recovery always decreased with increasing feed throughput or cycle time. In contrast, the purity always increased with increasing feed throughput with it tapering off at higher values of θ, while it went through a maximum with increasing cycle time. These trends were subtlety different for the cycles with LE (cycle Ia) and HE (cycle Ib) Eq steps. At larger θ and longer cycle times, HE Eq (cycle Ib) caused marginal improvement in purity while the recovery was slightly lower compared to that of LE Eq (cycle Ia). However, as θ decreased, cycle Ib exhibited sudden and very sharp decreases in purity for the 40 and 60 s cycle times. These trends were not nearly as drastic for cycle Ia. This comparison suggested LE Eq G
DOI: 10.1021/acs.iecr.6b02570 Ind. Eng. Chem. Res. XXXX, XXX, XXX−XXX
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Figure 7. Performance of the 3-bed 5-step PSA cycle schedule with the LE Eq step (Ia): effect of the valve coefficient (Cv) on the O2 purity and O2 recovery (a and c) and cycle step pressure history (b and d) during the Eq step (a and b) and during the CnD step (c and d) for a 10 SLPM feed flow rate and 60 s cycle time. The cycle step sequence is F-EqD-CnD-EqU-LPP.
longer saturated and essentially being underutilized near the light end. However, of course, the recovery kept increasing. The effect of cycle time on purity was due to the strictly kinetic nature of this O2:Ar separation process. At the beginning of the CnD step the gas leaving the bed mostly came from the void space inside the bed with its composition thus being close to the feed composition with little enrichment in O2. Gradually, the faster diffusing O2 started to desorb into the gas phase, thereby enriching the heavy product with O2. Toward the end of the CnD step, Ar now had enough time to desorb into the gas phase and contaminate the heavy product. Thus, the maximum observed in purity for the 60 s cycle time, relative to the 40 and 80 s cycle times, confirmed there was an optimum time where the step or cycle time-constant was in sync with the mass-transfer time-constant of Ar. Next, a valve coefficient (Cv) study was carried out with both pressure changing steps of cycle Ia (i.e., the Eq and CnD steps) to investigate their effects on the process performance and to obtain appropriate values to use with cycles II and III. The results are presented in Figure 7 for a wide range of Cv values in terms of purity and recovery and cycle step pressure history. For both cycle steps and by design, the largest Cv values investigated caused the pressure to decrease to final values in less than a second, and the smallest Cv values caused the pressure to decrease to final values just by the end of the 10 s step duration. For both cycle steps, the process performance was not sensitive to the Cv until smaller values were approached; the purity then began to decrease for both cycle steps, but more so for the CnD step, while the recovery
was better than HE Eq for this O2:Ar kinetic separation. Cycle Ia also resulted in the better overall process performance, as discussed earlier and shown in Table S5. These were interesting results because Jee at al.8 showed only that adding a HE Eq step improves the performance of this kind of separation, whereas this work showed a LE Eq step may even be better. This result also suggested using only LE Eq steps in the 4-bed 7-step cycles. With respect to trends for either cycle, increasing the cycle time resulted in a longer feed step, which in turn caused more loss of O2 into the light product during the feed step, which in turn caused lower recoveries in the heavy product. This was in spite of the fact that the durations of the other steps relative to the feed step did not change. This result suggested that at longer cycle times regeneration of the bed was insufficient to overcome the longer feed step. The effect of the feed throughput, i.e., increased feed flow rate, on recovery was now clearly caused by insufficient bed regeneration because more feed was fed to a bed with increasing θ without changing the relative durations of any of the regeneration steps. The effect of θ on purity was understood as follows. At larger θ the bed was essentially saturated at the feed conditions. This resulted in the maximum purity being attained at the largest θ but with the lowest recovery, as explained. Over a range of decreasing θ, the purity remained almost constant while the recovery increased significantly. This showed the system was losing less O2 into the light product but was still close to being saturated. When θ decreased even further, the purity exhibited a sharp decrease for all cycle times, which showed the bed was no H
DOI: 10.1021/acs.iecr.6b02570 Ind. Eng. Chem. Res. XXXX, XXX, XXX−XXX
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Industrial & Engineering Chemistry Research increased for the Eq step and decreased for the CnD step, both only slightly. For the CnD step, when using very low Cv values that corresponded to gradual changes in pressure, the driving force for desorption was lowered; thus, less O2 ended up in the heavy product, which led to lower purities and recoveries. The slower depressurization steps also allowed time for Ar to desorb and contaminate the heavy product, which also decreased purity. For the Eq step, lower Cv values resulted in lower purities and higher recoveries because now the O2 that did not desorb ended up staying in the bed and was recovered during the CnD step, but with Ar once again having more time to desorb and spoil the purity of the heavy product. These results suggested that Cv values in the insensitive range should be used; therefore, values of 0.1 and 0.2 were used for the Eq and CnD steps, respectively, in the simulations of the 4-bed 7-step cycles. Figure 8 shows the results for cycle II. The trends and the general behavior of this 4-bed 7-step PSA cycle were similar to
Figure 9. Performance of the 4-bed 7-step PSA cycle schedule III: effect of the CoD step end pressure on the O2 purity for different cycle times and at 10 SLPM feed flow rate. The cycle step sequence is FEqD-CoD-CnD-LEP-EqU-LPP.
step end pressure on the O2 purity for cycle III. A high feed flow (i.e., 10 SLPM) was chosen to ensure near saturation of the bed to achieve the maximum possible purity. CoD step end pressures purposely lower than the Eq step end pressures of cycle II were examined at three different cycle times. At all three cycle times, the purity exhibited a maximum with the CoD end pressure. The reason for this behavior was once again strictly due to the time-dependent nature of kinetic separation. When the pressure was lowered, both O2 and Ar desorbed and diffused into the gas phase. However, because O2 diffuses faster than Ar, the gas phase became enriched in O2 and the adsorbed phase became enriched in Ar. Therefore, if the CoD end pressure was too low, too much O2 was removed and recycled during the CoD step, leaving the bed with an adsorbed phase relatively enriched in Ar. The bed, which now contained less O2, produced a heavy product with lower purity during the CnD step. CoD end pressures of around 150 and 175 kPa resulted in the highest purities. In fact, the two best process performances of all the cycles in terms of purity and recovery were obtained from cycle III with a CoD end pressure of 175 kPa. These are runs 28 and 29 in Table S6, with purity >99.5 vol % and recovery >90%. Figure 10 displays the performance of cycle III for two different CoD step end pressures (i.e., 150 and 175 kPa). Notice in both panels the higher purities of O2 obtained from cycle III compared to those obtained from cycles Ia and II. In some cases over 99.8 vol % O2 was produced. This result clearly indicated that a further decrease in the bed pressure before the CnD step resulted in higher purities. However, during this CoD step, a significant amount of O2 was removed from the bed, and although it was recycled to the bed undergoing the LEP step, the recoveries still decreased rather markedly in comparison with cycle II when operated under similar conditions. Cycle III with a CoD end pressure of 175 kPa still provided the best overall process performance in terms of purity first and then recovery, as highlighted earlier. As mentioned earlier, there might be some N2 present in the O2 product of an equilibrium-based air separation PSA system. Two final studies were carried out, first to investigate the effect of N2 in the feed on the PSA process performance using two ternary feed mixtures, i.e., with O2:Ar:N2 ratios of 93.1:4.9:2.0 and 91.2:4.8:4.0 (vol %), and second to investigate the effect of
Figure 8. Performance of the 4-bed 7-step PSA cycle schedule II: effect of the feed throughput and cycle time on the O2 purity and O2 recovery. The cycle step sequence is F-EqD1-EqD2-CnD-EqU2EqU1-LPP.
those of the 3-bed 5-step PSA cycle (Ia), but with a noticeable difference in purity. Because the only difference between cycles Ia and II was an additional Eq step, the results showed that adding another LE Eq step had a significant effect on purity. Table S5 shows that the final pressure before the CnD step was in the range of 226 to 254.5 kPa for cycle II in comparison to 287 to 303.5 kPa for cycle Ia. Therefore, by including two LE Eq steps instead of one, the initial CnD step pressure was lowered, which removed and recycled more gas from the column void spaces that contained a great deal of Ar prior to producing the heavy product. This resulted in a heavy product more enriched in O2. In fact, as summarized earlier, cycle II exhibited the second best process performance based on purity first and recovery second. Cycle III utilizes a forced CoD/LEP step instead of the second LE Eq step. As mentioned before, this step was used in lieu of an Eq step because, unlike an Eq step, the end pressure of a CoD/LEP step was controllable. As mentioned earlier, the effect of this unique CoD/LEP step when also following an LE Eq step has also been studied by others for an equilibriumbased PSA separation.20 Figure 9 displays the effect of the CoD I
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Figure 10. Performance of the 4-bed 7-step PSA cycle schedule III: effect of the feed throughput and cycle time on the O2 purity and O2 recovery for CoD end pressures of (a) 175 and (b) 150 kPa. The cycle step sequence is F-EqD-CoD-CnD-LEP-EqU-LPP.
Figure 11. Performance of the 4-bed 7-step PSA cycle schedule III: (a) Effect of the feed throughput and feed composition on the O2 purity and O2 recovery for a 40 s cycle time and 175 kPa CoD end pressure. Binary feed contains an O2:Ar ratio of 95:5, and ternary 1 feed and ternary 2 feed contain 2 and 4 vol % N2, respectively, with the O2:Ar ratio fixed at 95:5. (b) Effect of the feed throughput and N2 replacing Ar in the feed on the O2 purity and O2 recovery for a 40 s cycle time and 175 kPa CoD end pressure. The ternary 2 feed contains O2:Ar:N2 ratios of 91.2:4.8:4.0, and binary 2 feed contains O2:Ar ratios of 91.2:8.8 in vol %. The cycle step sequence is F-EqD-CoD-CnD-LEPEqU-LPP.
replacing N2 with Ar in the feed, where the 4 vol % N2 in the 91.2:4.8:4.0 feed mixture was replaced with Ar to form a new binary mixture with an O2:Ar ratio of 91.2:8.8. In the first case, the O2:Ar ratio was kept constant and equal to the binary composition (95:5), which meant the O2 content in the feed decreased with increasing N2 content. In the second case, the O2 content in the feed was kept constant and the N2 content was simply replaced with Ar. These different feed mixtures were examined with cycle III at a 40 s cycle time and 175 kPa CoD end pressure because this cycle and those conditions produced the best results with the binary O2:Ar 95:5 feed mixture. The results are presented in Figure 11 in terms of O2 purity and O2 recovery for different feed throughputs. The results in Figure 11a show that for feeds containing 0, 2, and 4 vol % N2 a rather marked decrease in purity occurred with increasing N2 content (which also corresponded to decreasing O2 content), while there was essentially no effect on recovery. The results in Figure 11b show that when the O2 content in the feed was fixed, but with Ar replacing all the N2, there was essentially no effect on recovery and a very slight negative effect on purity. This subtle and intuitively unexpected effect on purity is explained below from a fundamental point of
view after explaining from the process performance point of view. From a process performance point of view, because there was essentially no effect of different levels of N2 and Ar at a fixed O2 content, together the results in Figure 11 indicated very clearly that the decrease in purity shown in Figure 11a was due to decreasing O2 levels in the feed. In fact, the results exhibited an almost proportionally negative effect on purity. This meant that the relative ratio of N2 to Ar in the feed did not matter. All that mattered to the process performance was the level of O2 in the feed, and it was critically important to obtaining a high-purity O2 product. From a fundamental point of view, because the N2 and Ar equilibrium properties are very similar (see Figure 4d) and because N2 diffuses about two times faster than Ar (see Table 1), intuitively, removing N2 from the feed should have caused J
DOI: 10.1021/acs.iecr.6b02570 Ind. Eng. Chem. Res. XXXX, XXX, XXX−XXX
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Figure 12. Effect of O2 on the binary mixed gas isotherms of Ar and O2, and N2 and O2 as predicted from the dual-process Langmuir model at 303 K for (a) 0, (b) 25, (c) 225, and (d) 500 kPa O2.
molecular sieve. The simulations were done with an experimentally validated dynamic adsorption process simulator (DAPS) developed in-house. The equilibrium isotherms for O2, N2, and Ar on Takeda CMS-T3A were obtained from the literature and fitted to the dual-process Langmuir model. A loading-independent isosteric heat of adsorption was obtained by fitting the same data to the Toth model. The mass-transfer rates of these adsorbates in CMS-T3A were measured in-house using a volumetric frequency response apparatus and correlated to a LDF model. Two binary (O2:Ar) and two ternary (O2:Ar:N2) feed mixtures were investigated to mimic the product produced from a typical O2 PSA system utilizing a molecular sieve zeolite. Besides feed composition, the effects of light end versus heavy end equalization, two sequential Eq steps versus Eq followed by forced cocurrent depressurization/light end pressurization steps, feed flow rate (i.e., feed throughput θ), cycle time, Eq step valve coefficient, and countercurrent depressurization step Cv on the PSA process performance were studied in terms of O2 purity (vol %), O2 recovery (%), and θ (LSTP/h/kg). The key process performance parameter was O2 purity because the PSA cycle had to produce ≥99.5 vol % at any O2 recovery or feed throughput to satisfy a constraint imposed by NASA, i.e., O2 containing less than 0.5 vol % Ar. The O2 recovery and feed throughput were considered to be equally important after O2 purity, because a higher O2 recovery means less power consumption while a higher θ means less volume and weight. In this study the best PSA cycle was arbitrarily defined as one that meets the O2 purity requirement and also exhibits the highest O2 recovery at any θ. In descending order of performance, the 4-bed 7-step PSA cycle with LE Eq and forced CoD/LEP (cycle III) resulted in an O2 purity of 99.5 vol % at an O2 recovery of 93.3% and θ = 455.7 LSTP/h/kg. This was followed by the 4-bed 7-step cycle
an increase in the O2 purity, but the opposite result was observed, i.e., the O2 purity was slightly but consistently lower without N2 in the feed over a wide range of θ, as shown in Figure 11b. The only plausible explanation for this result resided in a subtle mixed gas equilibrium isotherm effect. This notion was investigated by plotting binary mixed gas isotherms of Ar or N2 in the presence of different levels of O2. The results are displayed in Figure 12 based on predictions from the DPL model at 303 K. Figure 12a shows the two single-gas isotherms for Ar and N2 in the absence of O2; they were consistent with those shown in Figure 4d, in that Ar adsorbs slightly more than N2. However, when different levels of O2 were added to either Ar or N2 to create binary mixed-gas isotherms of Ar and O2, and N2 and O2, it became obvious that O2 suppressed N2 much more than it suppressed Ar. The effect became more pronounced at higher O2 partial pressures. Because of the dominant presence of O2 in the feed, the N2 loadings were much lower than the Ar loadings in a ternary mixture. This reduced the mass-transfer driving force for N2, which resulted in less desorption of N2 in comparison to Ar, thereby counteracting the effect of N2 having the larger mass-transfer coefficient. This interesting and subtle equilibrium effect was unexpected because the equilibrium isotherms are so similar. Nevertheless, it caused the O2 purity to be slightly lower when N2 was replaced by Ar in the feed, as shown in Figure 11b. This result was important from a fundamental point of view, but it was not as important from a process point of view, because the effect on the O2 purity was very small, as shown in Figure 11b.
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CONCLUSIONS A simulation study was undertaken to evaluate two 3-bed 5-step and two 4-bed 7-step continuous feed pressure swing adsorption cycle schedules systematically devised for producing high-purity O2 (≥99.5 vol %) at high recovery using a carbon K
DOI: 10.1021/acs.iecr.6b02570 Ind. Eng. Chem. Res. XXXX, XXX, XXX−XXX
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with two LE Eqs (cycle II), which provided an O2 purity of 99.6 vol % at an O2 recovery of 85.7% and θ = 471.4 LSTP/h/kg. These two performances were similar except for recovery. In contrast, the 3-bed 5-step cycle with LE Eq (cycle Ia) provided an O2 purity of 99.5 vol % at a θ = 1089.5 LSTP/h/kg but with an O2 recovery of only 77.4%. The worst performing PSA cycle was the 3-bed 5-step cycle with HE Eq (cycle Ib), with it resulting in an O2 purity of 99.5 vol % at an O2 recovery of 68.7% and a θ of 628.6 LSTP/h/kg. The two 3-bed 5-step PSA cycle schedules exhibited the highest feed throughputs, but the corresponding recoveries were low. The additional Eq or CoD/ LEP cycle steps that required an additional bed to be added to the 3-bed PSA cycle schedule significantly improved the O2 recovery at the expense of decreasing the feed throughput. In general, it was easy to obtain an O2 purity >99.5 vol % at an O2 recovery >50%; however, achieving an O2 purity >99.5 vol % at an O2 recovery >90% was challenging but doable. The 4-bed 7step PSA cycle with LE Eq and forced CoD/LEP (cycle III) achieved it just twice from a total of 102 case studies with all PSA cycles. There were no effects of the two different Cv values (i.e., those used during the Eq and CnD steps) on the process performance until smaller values were used, and then the effects were only marginally unfavorable on O2 purity. All four PSA cycles were very sensitive to both tc and θ, with O2 purity decreasing and O2 recovery increasing with decreasing θ, and O2 recovery decreasing and O2 purity exhibiting a maximum with increasing tc. For cycle III, the O2 purity exhibited a maximum with a change in the CoD end pressure, with more pronounced negative effects on the O2 purity at longer tc. Replacing Ar with N2 in the feed at a fixed O2 content had essentially no effect on O2 recovery and a slightly negative effect on O2 purity, while decreasing the level of O2 in the feed at a fixed O2:Ar ratio had an almost proportionally negative effect on the O2 purity and no effect on O2 recovery. From a process performance point of view, this meant the relative ratio of N2 to Ar in the feed did not matter; all that mattered was the level of O2 in the feed, and it was critically important to achieving a high-purity O2 product. From a fundamental point of view, a very subtle and counterintuitive effect was observed, wherein higher levels of N2 relative to Ar in the feed resulted in this slightly higher O2 purity even though N2 diffuses twice as fast as Ar. This was explained by a counteracting and surprising equilibrium effect due to Ar adsorbing just slightly more than N2.
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ACKNOWLEDGMENTS
Continued financial support provided over many years by both the NASA Marshall Space Flight Center and the Separations Research Program at UT-Austin is greatly appreciated.
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ASSOCIATED CONTENT
S Supporting Information *
The Supporting Information is available free of charge on the ACS Publications website at DOI: 10.1021/acs.iecr.6b02570. Discussion of literature PSA cycles, initial and boundary conditions of mathematical model, mass-transfer coefficient determination from volumetric frequency response data, and tabulated DAPS results (PDF)
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AUTHOR INFORMATION
Corresponding Author
*E-mail:
[email protected]. Notes
The authors declare no competing financial interest. L
NOMENCLATURE bi = Toth model parameter for component i in eq 14, kPa−1 bo,i = Toth model parameter for component i in eq 15, kPa−1 B 1,i = dual-process Langmuir model parameter for component i in eq 6, K−1 B 2,i = dual-process Langmuir model parameter for component i in eq 7, K−1 b 1,i = dual-process Langmuir model parameter for component i in eq 6, kPa−1 b 2,i = dual-process Langmuir model parameter for component i in eq 7, kPa−1 0 b 1,i = dual-process Langmuir model parameter for component i in eq 6, kPa−1 0 b 2,i = dual-process Langmuir model parameter for component i in eq 7, kPa−1 Cp,g,i = gas-phase heat capacity of component i, kJ mol−1 K−1 Cp,a,i = adsorbed phase heat capacity of component i, kJ mol−1 K−1 Cp,g = gas-phase heat capacity, kJ mol−1 K−1 Cp,p = adsorbent particle heat capacity, kJ mol−1 K−1 CT = total molar concentration, mol m−3 Cv = valve coefficient, dimensionless Ei = mass-transfer activation energy for component i in eq 4, K−1 F = molar flow rate through the valve, L(STP)/min hw = overall heat-transfer coefficient, kW m−2 K−1 ΔHi = isosteric heat of adsorption of component i from eq 15, kJ mol−1 ki = mass-transfer coefficient of component i in eq 3, s−1 k0,i = mass-transfer pre-exponential factor for component i in eq 4, s−1 L = column length, m Mg = gas-phase average molecular weight, kg mol−1 N = number of components P = pressure, kPa P0 = pressure outside the valve, kPa qi = adsorbed phase loading of component i, mol kg−1 q*i = adsorbed phase equilibrium loading of component i, mol kg−1 qsi = Toth model parameter for component i in eq 14, mol kg−1 s q 1,i = dual-process Langmuir model parameter for component i in eq 5, mol kg−1 s q 2,i = dual-process Langmuir model parameter for component i in eq 5, mol kg−1 R = universal gas constant, kPa m3 mol−1 K−1 rb,i = column internal radius, m rp = adsorbent particle radius, m Sg = gas-phase specific gravity relative to air at 1 atm and 21.45 °C t = time, s ti = Toth model parameter for component i in eq 14 T = temperature, K T0 = ambient temperature, K v = interstitial velocity, m s−1 yi = mole fraction of component i z = column axial coordinate, m DOI: 10.1021/acs.iecr.6b02570 Ind. Eng. Chem. Res. XXXX, XXX, XXX−XXX
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Industrial & Engineering Chemistry Research Greek Symbols
Recovery from Flue Gas By a Novel Vacuum Swing Adsorption Cycle. Comput. Chem. Eng. 2011, 35, 1010−1019. (16) Brown, P. N.; Hindmarsh, A. C.; Petzold, L. R. Using Krylov Methods in the Solution of Large-Scale Differential-Algebraic Systems. SIAM J. Stat. Comp. 1994, 15, 1467−1488. (17) Giesy, T. J.; LeVan, M. D. Mass Transfer Rates of Oxygen, Nitrogen, and Argon in Carbon Molecular Sieves Determined by Pressure-Swing Frequency Response. Chem. Eng. Sci. 2013, 90, 250− 257. (18) Ritter, J. A.; Bhadra, S. J.; Ebner, A. D. On the Use of the Dual Process Langmuir Model for Correlating Unary and Predicting Mixed Gas Adsorption Equilibria. Langmuir 2011, 27, 4700−4712. (19) Bhadra, S. J.; Ebner, A. D.; Ritter, J. A. On the Use of the Dual Process Langmuir Model for Predicting Unary and Binary Isosteric Heats of Adsorption. Langmuir 2012, 28, 6935−6941. (20) Bae, Y. S.; Lee, C. H. Sorption Kinetics of Eight Gases on a Carbon Molecular Sieve at Elevated Pressure. Carbon 2005, 43, 95− 107. (21) Hossain, M. I. Volume Swing Frequency Response Method for Determining Mass Transfer Mechanisms in Microporous Adsorbents; Ph.D. Dissertation, University of South Carolina, Columbia, SC, 2014. http://scholarcommons.sc.edu/etd/2585. (22) Yasuda, Y. Frequency Response Method for Study of the Kinetic Behavior of Gas-Surface System. 1. Theoretical Treatment. J. Phys. Chem. 1976, 80, 1867−1869.
εb = column porosity εp = adsorbent particle porosity ρa,i = adsorbed phase density of component i, kg m−3 ρp = adsorbent particle density, kg m−3 μg = gas-phase viscosity, Pa s Cycle Step Acronyms
CnD = countercurrent depressurization CoD = cocurrent depressurization CoLR = concurrent light reflux Eq = equalization EqD = equalization down EqU = equalization up F = feed FP = feed pressurization HE = heavy end HR = heavy reflux I = idle LE = light end LEP = light end pressurization LPP = light product pressurization LR = light reflux
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REFERENCES
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DOI: 10.1021/acs.iecr.6b02570 Ind. Eng. Chem. Res. XXXX, XXX, XXX−XXX
